7 TEPELNÉ ZPRACOVÁNÍ NÁSTROJŮ V PROSTŘEDÍ INERTNÍCH PLYNŮ HEAT TREATMENT OF TOOLS IN NOBLE GAS ATMOSPHERE Mladen Stupnišek, Boţidar Matijević University of Zagreb, Faculty of Mechanical Engineering and Naval Architecture, Ivana Lučića 5, Zagreb, Croatia, ABSTRACT V tomto příspěvku je popsán vývoj zařízení pro kalení nástrojů. Je dobře známo, ņe s rostoucí ochlazovací intenzitou během kalení klesá výskyt precipitace karbidů na hranicích zrn. V důsledku toho se zhorńují dynamické vlastnosti nástrojů. S ohledem na tuto skutečnost se vývoj vakuových pecí vydává dvěma směry. Jeden směr je zaloņen na zvyńování tlaku plynu a hledání vhodnějńích plynů pro kalení v přetlaku plynu. Jiný směr spočívá ve vývoji zařízení pro kalení nástrojů zaloņeném na ohřevu v atmosféře inertních plynů (argon nebo helium) s následným zakalením v termální lázni, čímņ je moņné vyrobit technologicky zdokonalené nástroje s niņńími náklady neņ mají vakuové technologie. Klíčová slova: nástrojové oceli, kalení, precipitace karbidů, inertní plyn, rychlost kalení The development of equipment for tool hardening is described. It is a well known fact that an increase in cooling intensity during quenching decreases the occurrence of carbide precipitation on grain boundaries. Consequently, dynamical properties of tools in exploitation are deteriorated. With that in mind, the development of vacuum furnaces is oriented in two directions. One direction is based on an increase in gas pressure and a search for better suited gasses for high pressure gas quenching. The other direction is the development of tool hardening equipment based on heating in a noble gas atmosphere (argon or helium) followed by quenching in a proprietary thermal bath basin, thus making the tools produced more technologically advanced and cost-effective in comparison to those produced by the vacuum furnace technology. Key words: tool steels, hardening, carbide precipitation, noble gas, quenching rate 1. EQUIPMENT FOR TOOL HARDENING Widely used atmosphere integrated quench furnaces are rarely applied for tool hardening due to the inadequate protection of steel surface, or, on the other hand, due to the reaction of active gas components with carbon in the steel surface, which may cause decarburizing or carburizing. Even nitrogen from the gas atmosphere may change the chemical composition of the steel surface by absorption in austenite, which is then stabilized. Consequently, an increased amount of retained austenite is present in the hardened steel. The application of conventional atmosphere furnaces is bad for the environment because they use natural gas sources and emit harmful gas components, such as carbon dioxide and nitrogen oxides. Vacuum furnaces are mostly used for tool hardening due to the protection of steel surface during heating, but the limited cooling intensity during gas quenching is the main

8 disadvantage of conventional vacuum furnaces. High alloyed tool steels have high hardenability, as shown in Fig. 1, [1]. Due to a high amount of carbon in many cases, the precipitation of pre-eutectoid carbides results in decreased toughness and thermal fatigue resistance of tools, [2]. Fig. 1. CTT diagrams of selected tool steels [1] Carbide precipitation Martensite Fig. 2. Carbides precipitated on grain boundaries [2] Low carbon martensite near grain boundaries is the cause of lower resistance to tempering, which consequently decreases hardness during the high temperature application. Fig. 3. Influence of quenching intensity on the softening of hot work tool steel under the surface during die casting [3]

9 In the past few years, the quenching intensity of vacuum furnaces has been notably increased by the use of high pressure gas quenching or the integration of oil bath quenching. Figure 4 shows the influence of the kind and the pressure of gas on heat transfer in comparison to oil bath quenching. Fig. 4. Comparison of high pressure gas quenching and oil bath quenching [4] Fig. 5. Vacuum furnace with oil bath quenching [5]

10 Vacuum furnaces have been developed in two main directions. One direction is based on an increase in gas pressure and the substitution of nitrogen gas with helium or hydrogen, whereas the other direction incorporates vacuum oil quenching technology [5]. 2. A NEW TYPE OF ATMOSPHERE FURNACE FOR TOOL HARDENING In common integrated quench furnaces the active gas atmospheres that are used cannot fully protect the steel surface and decarburization or carburization take place on the surface of tools. Also, nitrogen is not absolutely inert because, at high temperatures, it absorbs the stabilizing austenite into the steel surface. Consequently, the amount of retained austenite in hardened tools increases. The retained austenite decreases hardness and compressive stresses in the surface layer. Only noble gases are absolutely inert to steel but their application in common quench furnaces is not adequate due to high consumption of expensive noble gases. To make the application of noble gases (argon and helium) possible, a new design of integrated quench furnace has been developed in order to enable low consumption of expensive gases. The design of the new type of furnace is based on the fact that densities of argon and helium are significantly different from the density of the air, which enables simple atmosphere changes based on the gravity principle to be made without evacuation. Argon is 40 % heavier than air and helium is 7 times lighter than air. Fig. 6. Functional schema of integral quench furnace for thermal bath hardening from the helium gas atmosphere [6]

11 The new integral quenching furnace is a vertically arranged bell-type furnace with an enclosed vestibule sealed on a quenching basin with a lift for vertical transport of charge in a closed gas system [6]. Helium is placed in the upper part of the furnace. During heating, it expands down through the vestibule to the expanding vessel where it resides until the start of cooling cycle when it comes back into the vestibule. The vertical transport of charge from the furnace through the vestibule into the quenching basin is carried out without any contact with air and the steel surface remains perfectly bright. The whole hardening process, i.e. heating and quenching, can be driven by the data obtained from two thermocouples built in the test block, measuring the surface and the core temperature. For a quenchant, it is possible to choose hot oil bath or salt thermal bath, or even water because its vapour cannot damage the helium gas atmosphere. Gas atmosphere pit-type furnaces are suitable for hardening in argon because argon is heavier than air and the flow direction is from the bottom to the top of the retort. Furnaces for tempering are designed by applying the same principles as for gassing retorts. 3. CONCLUSION The developed integral quench furnace enables hardening in liquid media without any influence on the steel surface. The quenching intensity of liquid media is very high and it is comparable with that of high pressure gas quenching in vacuum furnaces using helium or hydrogen at a pressure of more than 20 bars. Besides technological advantages, the developed equipment is more economically justified by lower investment and exploitation costs. 4. LITERATURE [1] Prospect of K340, W300 and S600, [2] A. Dietrich, E. Haberling, K. Rasche, I. Schruff: Alloy Optimization of Hot-Work Tool Steel. Thyssen Edelstahl technische Berichte 1990, [3] D. Magnaca: Heat treatment of hot-work tool steels. 7 th Tooling conference, Politechnico di Torino, 2-5 May [4] P. Kula, R. Pietrasik, K. Dybowski, R. Atraszkiewicz, E. Stanczyk- Wolowiec, M. Korecki, J. Olejnik: New technological pathways for universal vacuum furnaces, 18 th IFHTSE Congress, Rio de Janeiro, Brazil, 26 th to 30 th June 2010., , CD [5] Prospect of Fours BMI, Vacuum furnace for oil quenching, [6] M. Stupnińek: Heat treatment in an environmentally-friendly gas atmosphere, 18 th IFHTSE Congress, Rio de Janeiro, Brazil, 26 th to 30 th June 2010., , CD

12 SAMOMAZNÉ TENKÉ POVLAKY PRO NÁSTROJOVÉ OCELI SELF-LUBRICATING THIN FILMS FOR TOOL STEELS Peter Jurči a, Stanislav Krum b a MtF STU Trnava, Paulínská 16, Trnava, Slovac Republic b ČVUT v Praze, Fakulta strojní, Karlovo nám. 13, Praha 2, Czech Republic ABSTRACT Vzorky vyrobené z oceli pro práci za studena Vanadis 6 byly vyrobeny třískovým opracováním, brouńeny a tepelně zpracovány standardním reņimem pro daný materiál. Nakonec byly vzorky leńtěny do zrcadlového lesku. V tomto stavu byl materiál povlakován povlaky CrAgN o koncentracích stříbra 3 hm.% a 15 hm.%. Mikrostrukturní analýzou vylo zjińtěno, ņe přísada 3 hm.%ag nemá významný vliv na mechanismus růstu povlaku. Na druhé straně přísada 15 hm.%ag má na růst povlaku významný vliv. Povlak s 3 hm.% Ag má vynikající adhezi na základní materiál. Oba typy povlaků mají výborné tribologické vlastnosti při tření s protikusy z Al2O3 a litého cínového bronzu CuSn6. Specimens made from Vanadis 6 cold work tool steel were machined, ground, heat processed by standard regime and finally mirror polished. After that, they were layered with CrAgN. The Ag-content in the layers was chosen to 3 wt% and 15 wt% respectively. Microstructural analysis revealed that the addition of 3 wt%ag did not influence the growth manner of the films but the addition of 15 wt%ag has made considerable changes in the film growth. The layer with 3 wt%ag had excellent adhesion on the steel substrate. Both films have also superior tribological properties against hard material (alumina) as well as against soft counterpart (CuSn6 as-cast bronze) 1. INTRODUCTION Thin CrN-based films have been used in variety of industrial applications like copper machining, aluminium die casting and forming and wood processing [1-7]. However, some of tribological properties of these films cannot be changed in a sufficiently wide range since they are given by the nature of the film compound itself. Therefore, the effect of selflubrication gained a great scientific importance in last few years. The main idea to develop self-lubricating and multi-purpose coatings was based upon the fact that commercially available lubricants (sulphides, oxides, graphite) exhibit considerable shortcomings and can not be used effectively in tooling applications. Soft noble metals, on the other hand, posse stable chemical behaviour and can exhibit self-lubricating properties due to their low shear strength. Self-lubricating effect is based on incorporation of small amount of noble metals, mostly silver, into the basic CrN-film. Silver is completely insoluble in CrN and forms nanoparticles in basic CrN-compound. These particles are stable up to high temperatures, have low hardness and shear strength and do not behave as abrasives. Siler containing transition metal nitrides films have been extensively studied in recent years. Various authors established that the addition of small amount of silver into ZrN, YSZ and CrN increases the H 3 /E 2 ratio [8], wear resistance [9] and significantly reduces the friction coefficient [10-13].

13 CrAgN coatings deposited on various substrates, but excepting ledeburitic tool steels, were studied by several authors [11-13]. They established alterations in the CrN-layer growth orientation with increased Ag-content [12]. The friction coefficient of the layers with 22 at.% Ag, measured for the 100Cr6-steel counterpart at 600 o C was reduced from 0.64 (pure CrN) to 0.47 (CrN+22 at.% Ag) [11]. As reported elsewhere [13], silver diffuses to the surface at high temperature, forms lubricious grains there, which gives the principal explanation of superior tribological properties of the CrAgN-films. Current paper deals with the development of adaptive nanocomposite CrAgN coatings on the Vanadis 6 Cr-V ledeburitic tool steel. It describes and discusses the basic coating characteristics like wear resistance, friction coefficient, Young s modulus, as a function of the silver content and deposition temperature. 2. EXPERIMENTAL 2.1. Material and processing The experimental samples were made from the ledeburitic steel Vanadis 6 with nominally 2.1 %C, 1.0 %Si, 0.4 %Mn, 6.8 %Cr, 1.5%Mo, 5.4 %V and Fe as balance. After rough machining procedure to the semi-final dimensions (plates 55 x 10 x 1 mm), they were subjected to standard heat treatment procedure. After that the samples were fine ground and polished with the diamond suspension up to the mirror finish. The CrN- and CrN/Ag - coatings were deposited in a magnetron sputter deposition system, in a pulse regime with a frequency of 40 khz. Two targets, opposite positioned, were used: One made from pure chromium (99.9%Cr) and the second made from silver of 99.98% of purity). The cathode output power was 5.8 kw on the chromium cathode. On the silver cathode, the output powers were 0.1 and 0.45 kw in order to produce the silver contents in the coating of 3 wt% and 15 wt%, respectively. The processes were carried out in a low pressure atmosphere (0.15 mbar), containing the nitrogen and the argon, in a ratio of 1:4.5. The substrates were placed between the targets on rotating holders, with a rotation speed of 3 rpm. Just prior the deposition, the substrates were sputter cleaned in an argon low pressure atmosphere for 15 min. The substrate temperature was 250 o C for the cleaning. For the deposition, the temperature was increased to 500 o C using an internal wall heating. Negative substrate bias of 200 V was used for the sputter cleaning and that of 100 V for the deposition. The total deposition time was 6 hours Investigation methods Microstructural analysis has been carried out on fracture surfaces of coated samples, on the field emission scanning electron microscope JEOL JSM-7600F. Coated specimens were immersed into liquid nitrogen and broken down before the analysis. The nanohardness and the Young s modulus (E) of the coatings were determined using the instrumented nanoindentation test under a normal load of 20 mn, at a NanoTest (Micro Materials Ltd) nanohardness tester equipped with a Berkovich indenter. For both coatings, ten measurements were made and the mean value and the standard deviation were calculated. The adhesion of the coatings on the substrate has been evaluated using a CSM Revetest scratch-tester. The scratches were made under progressive increasingly load from 1 N to 100 N, with a loading rate of 50 N/min. Standard Rockwell diamond indenter with a tip radius of 200 mm was used. Five measurements were made on each specimen and the mean value of adhesion, represented by the L c1 and L c2 critical loads, respectively, has been calculated. The critical loads were determined by the recording of the signal of acoustic emission as well as by the viewing of the scratches on the light micrographs. The L c1 critical load corresponded to the occurrence of first inhomogenities in the coating and the L c2 critical load was determined as load when 50% of the coating was removed from the substrate.

14 Tribological properties of the coatings were measured using the CSM Pin-on-disc tribometer, at ambient and elevated temperatures, up to 500 o C. For the testing, the balls with a 6 mm in diameter, made from sintered alumina and CuSn6 bronze have been used. No external lubricant was added during the measurements. The normal loading used for the investigations was 1 N. For each measurement, the number of cycles was 5100, e.g. at the sliding radius of 5 mm, the total sliding distance was 100 m. 3. RESULTS AND DISCUSSION The microstructure of the substrate material after the heat treatment is in Fig. 1. The material consists of the matrix, formed with tempered martensite and fine carbides, uniformly distributed throughout the matrix. As established recently [14] that it is mainly the M 7 C 3 - phase that underwent the dissolution in the austenite during the heat processing. This results in the saturation of the austenite with carbon and alloying elements, which leads to high hardness of as-heat treated material. Other part of the M 7 C 3 -carbides, and almost complete amount of MC-phase remained undissolved. After the heat treatment, the average hardness of the material was 724 HV m Figure 1. Microstructure of PM ledeburitic steel Vanadis 6 substrate after heat treatment a 3 m 3 m Figure 2. SEM micrographs showing the microstructure of CrAg3N (a) and CrAg15N (b) films developed on the Vanadis 6 steel substrate The thickness of the film with 3%Ag addition was 4.2 m, Fig 2a. Previous paper [27] was devoted to the analysis of CrAgN films prepared at the temperature of 260 o C (corresponding to heating up due to ion bombardment only). It has been established that the b

15 films formed at low temperature have a similar thickness, e.g. higher deposition temperature does not influence the growth rate significantly at a given Ag addition. On the other hand, the film with an addition of 15 wt% Ag has grown much faster and its total thickness was established to be 6.3 m, Fig. 2b. The film with 3 wt% Ag addition grew in a columnar manner, with well visible individual crystallites, Fig. 2a. There is no difference in a growth manner between this film and that formed at 260 o C, as reported previously [15]. The situation in the case of the film with an addition of 15 wt% Ag is clearly different individual columnar crystals are not visible but separate Ag-particles can be shown on the SEM micrograph, Fig. 2b. a 1.5 m b Figure 3. SEM micrograph showing the surface microstructure of CrAg15N film (a) and corresponding EDS-map of silver (b) SEM micrograph in Fig 3a, made from the surface in the BE-detection regime, and corresponding EDS mapping of Ag, Fig 3b show that silver forms individual grains on the surface at higher concentration. Previous paper [15] did not documented this fact mainly due to the fact that lower temperature was used for the deposition and the silver concentration was too low for the formation of individual crystals of sufficiently large size to be detected by SEM. In the paper published previously [15] it was concluded that the addition of 3 wt% Ag induced only very slight coating hardness decrease. The nanohardness of pure CrN was ± 1.49 GPa and that of CrAg3N ± 1.44 GPa. Current results confirmed that small Agaddition does not result in significant changes in coating hardness the hardness of 3 wt% Ag containing film deposited at 500 o C was ± 1.83 GPa, Table 1. The addition of 15 wt%ag, on the contrary, led to substantial hardness reduction it was only ± 0.61 GPa. Similar effect of the silver addition on the coating hardness has also been reported by Yao et al. [16] for magnetron sputtered nanocomposite coatings with various silver additions. An explanation can be given by the fact that Ag is very soft in nature and the Ag-particles embedded in CrN make a softening of the coating. Also the size of Ag-particles should be probably considered the larger the Ag-grains are the more considerable softening of the film. The Young s modulus of pure CrN and CrAg3N deposited at 260 o C were of about 240 GPa [15]. Also, small Ag addition in the film formed at 500 o C did not change the Young modulus, see also Table 1. But, higher silver amount incorporated into the basic CrN has a negative effect on E. The investigations published by Aouadi et al. [8], where decreased Young s modulus with increased Ag-content in YSZ-based coatings has been established, has thus been confirmed.

16 Table 1. Mechanical properties of the layers deposited at 500 o C Coating Hardness [GPa] Young s modulus [GPa] CrAg3N ± ± 17 CrAg15N ± ± 6 For the film with 3 wt%ag addition, the first symptoms of coating damage occurred at the average loading of around 47 N (L c1 ). Coating damage begins with an appearance of semicircular tensile cracks, Fig. 4a. The total failure of the CrAg3N - film is in Fig. 4b. It is typical by the occurrence of many parallel cracks in the scratch, where of about 50% of the coating is removed from the substrate. Typical load when this symptom occurred was N (L c2 ). a b c Figure 4. Light micrographs showing the failures after scratch test: a CrAg3N, L c1, b CrAg3N, L c2, c CrAg15N, L c1, d CrAg15N, L c2 Relatively softer 15 wt% of Ag containing coating fails also by the manner of semicircular tensile cracks, but the distance between the cracks is much larger than that in the CrAg3N-film, Fig. 4c. The critical load at which these symptoms firstly occurred was very low it ranged around 6.4 N, Table 2. Figure 4d shows the total failure of the CrAg15N coating. Here, higher distance between adjacent cracks is also clearly evident. One can assume that softer CrAg15N - coating can store a higher amount of plastic energy before failure. This seems to be logical since silver is soft in nature, forms individual particles in the coating and can make the coating more resistant to the brittle failure. Total failure of the d

17 coating with the 15 %Ag addition became at much lower load than that of 3 wt%ag containing film, Table 2. Compared the obtained results to those recorded from the coatings developed at 260 o C it seems that elevated deposition temperature (500 o C) has favourable effect to the CrAg3N adhesion. For the coating with 15 wt%ag, weak adhesion has been recorded. It seems that high silver content makes the coating too soft and very sensitive to the failure at higher loading. Table 2. Critical loads for defined degree of coatings failure Coating L c1 [N] L c2 [N] CrAg3N 46.9 ± ± 8.4 CrAg15N 6.4 ± ± 6.3 Table 3. Friction coefficients of coatings against two different counterpart s materials Testing temperature CrAg3N CrAg15N [ o C]/coating Al 2 O 3 CuSn6 Al 2 O 3 CuSn6 Room temperature Table 3 shows results of the measurements of friction coefficients. In all the cases, increased testing temperature led generally to decrease of the friction coefficient. At a room temperature, however, the friction coefficient against alumina was practically the same and, in addition, it does not exhibit any differences compared to pure CrN [15]. Therefore, our previous findings [15] were confirmed, e.g. no positive effect of the Ag-addition can be expected in low temperature applications. At higher operation temperature, on the contrary, there was positive effect of the silver addition found. This effect is slightly more evident for the 15 wt %Ag containing films. The friction coefficient against bronze was lower for both coatings at a room temperature. There is, unfortunately, no direct comparison to the pure CrN available. At higher temperature, decrease of friction coefficient has been recorded, which is more significant for the film with higher Ag-content. Practical application of the coating with higher Ag-content is, however, doubtless. On the one side higher Ag-content induces an improvement of friction characteristics for both types of counterparts. On the other side, however, the film with high Ag-content had poor adhesion on the substrate. In addition, increasing Ag-content makes the film deposition process more expensive. CONCLUSIONS Investigations of magnetron sputtered CrN-films with various Ag-additions have brought the following findings: The films with 3%Ag and 15%Ag formed at 500 o C had a thickness of 4.2 m and 6.3 m, respectively. Compared to the results obtained on the films developed at 260 o C it can be concluded that while the deposition temperature did not affect the final coating thickness, higher Ag-content led to greater thickness of the film.

37 quenchants may be used in different conditions (bath temperatures and agitation rates), thus contributing to the immense number of possible combinations. There are also different quenching techniques: direct immersion quenching; intensive quenching; interrupted quenching; delayed quenching; martempering; austempering; spray quenching. Yet there is no generally recognized method and technique for the measurement, recording and comparison the relative cooling intensities of different quenchants. From the other side computer modeling and simulation enable today scientifically based prediction of quenchhardness, microstructure, stresses and distortion in quenched workpieces. An indispensable prerequisite for any computer model is adequate heat transfer data for the entire duration of the quenching process, for actual quenching medium and quenching conditions. Currently, for measuring and recording the cooling intensity of liquid quenchants, the tests used are predominantly laboratory type (ISO 9950 and ASTM D 6200 for oils, and ASTM D 6482 and ASTM D 6549 for polymer-solutions), using a small specimen of 12.5 mm diameter 60 mm. The results of these tests (temperature/time and temperature/cooling rate plots see Fig. 1) are not directly applicable to real engineering components, because they do not provide heat transfer data for the complete quenching process of a real workpiece, and because the workshop conditions at quenching widely differ from those in laboratory tests. There is need to have a database of cooling intensities, also in workshop conditions, for different kinds of liquid quenchants, as a tool for designers and workshop engineers, available worldwide. Fig. 1. Typical a) temperature/time and b) temperature/cooling rate plots for test specimen 12.5 mm diameter 60 mm quenched in oil, according to ISO 9950 Industrial quenching oils Determination of cooling characteristics Laboratory test method a) b) The absence of adequate heat transfer data results in many cases of less than optimal selection of quenchant and quenching conditions. This leads to poor quality of manufactured workpieces, rework or even scrap. Hence, millions of dollars could be saved in manufacture, if designers and engineers in the workplace had adequate comparative data for cooling intensities of different quenching media with specified conditions. Given such a database, as a tool for computer programmers, computer modeling and prediction of quench-hardness, microstructure, stresses and distortion can become a normal practice when quenching real engineering components of complex geometry. The foreseen project for development of the database is two phase: Phase 1: Compiling the experimental results from different investigators, and establishing the database. Phase 2: Further development of computer models (3-D) for engineering components with complex geometry, using the results from Phase 1. Compiling the database, a reasonable number of selected quenchants will be included, every of them with specified conditions as they are normally used. In [8] Kobasko suggests: To use small silver probe for evaluation of critical heat flux densities [1]. To calculate the heat transfer coefficient based on testing by the Liscic/Petrofer probe and solving the inverse problem [2].

38 To apply the noise control system for investigation of transient boiling processes [3]. Taking into account very complicated cooling mechanisms from the one side, and the complex shape of the quenched workpiece (having thin and thick cross-sections, many holes and non-symmetric cavities), from the other side one has to be aware how difficult task is to provide real heat transfer data at quenching in workshop practice. This situation becomes even more complicated because, as it is known, the heat flux and consequently the heat transfer coefficient (HTC) vary around the workpiece's surface, which in some cases means different duration of the vapour-film phase at different points, and this leads to distortion. Because we cannot measure the temperature at so many different points on the surface, the question is: What is to be done to obtain real heat transfer data using an appropriate temperature measuring and recording method? The possible solution should incorporate: a) The probe itself should be of similar mass and shape as the workpiece to be quenched. This means that for cylindrical workpieces a cylindrical probe, for plate like workpieces a plate like probe, and for ring like workpieces a ring like probe, should be used. b) To calculate the HTC in different points on the surface of workpieces having complex geometry, an new advanced 3-D inverse heat conduction model, combined with local heat flux densities for different points, would probably be a possible solution. For the most frequently used liquid quenchants the database will give the following information: Description of chemical and physical properties; Specification of test conditions; Incorporate own testing methods; Laboratory test ISO 9950 with resulting cooling and cooling rate curves, and heat transfer coefficient as function of surface temperature, calculated using a uniformly predetermined mathematical procedure; Workshop test using the Liscic/Petrofer probe of 50 mm dia. x 200 mm with three resulting cooling curves and calculated heat transfer coefficient as function of surface temperature and of time, respectively; Critical heat flux densities determined by e.g. Japanese New Silver Probe, according to JIS method B; Noise test to determine transition from film to nucleate boiling processes. By applying the newest theory and methods, the foreseen database would supply comprehensive information about cooling characteristics of different quenchants, some of them not yet widely known, and enable an accurate comparison among them. Development of contemporary quenching technologies and testing methods, is taking place at many institutions/companies. Some of them from Europe, USA and Japan are widely known. From the other side we do not know enough about relevant developments in fast growing economies Brasil, Russia, India, China, which actually constitute the major part of potential database users. Because nowadays the industrial production is often transferred from one country to another, the database should be global, i.e. available to everyone in the world. During the 19 th Congress of the International Federation for Heat Treatment and Surface Engineering (IFHTSE) held on October 2011 in Glasgow, U.K. the Liquid Quenchants Database Project has been officially launched, and Dr. Imre Felde from Hungary was named as its leader. The immediate activity is to establish the project Consortium and its Core Group which will be responsible for: General operational management; Detailed planning; Communications within IFHTSE, public information dissemination and progress reporting.

39 2. THE LUMPED-HEAT-CAPACITY METHOD FOR CALCULATION OF THE HEAT TRANSFER COEFFICIENT Fundamental concept of this simple method, according to [4] is the following: If the probe temperature is uniform, the heat loss from the probe Q is equal to the decrease in the internal energy of the probe dt (1) p p l Q ha T T c V d t where h is the heat transfer coefficient on the probe surface, A is the surface area of the probe, T p the probe temperature, T l the quenchant temperature, c the specific heat of the probe material, ρ the specific density of the probe material, V the volume of the probe, t time, and dt p /dt the cooling rate of the probe. If the quenchant temperature around the probe T l is uniform, the next relation is derived from equation (1): V dtp q h Tp Tl c (2) A d t where q is the heat flux on the probe surface, and V dtp dt h c A Tp T (3) l The heat transfer coefficient can be directly calculated from the cooling rate dt p /dt, therefore the preciseness of q and h values depends on the accuracy of the cooling rate calculated from measured cooling curve data. The Lumped-Heat-Capacity Method can be used only if we can justify the assumption of uniform probe temperature during the cooling process. Temperature distribution in a probe depends on the thermal conductivity of the probe material and the heat transfer from the surface of the probe to the quenchant. In general, the smaller probe size and the higher the thermal conductivity of the probe material, the more realistic the assumption of uniform temperature of the probe. According to Kobasko it is assumed that a temperature field in a section of the silver probe is uniform, if the Biot number B i = h R/λ < 0.2, where R is the radius of the probe and λ is thermal conductivity (for silver λ is 16 times higher than for stainless steel or for Inconel). Fig. 2a shows the Japanese New Silver Probe of 10 mm diameter and 30 mm length, according to JIS 2242-method B; Fig. 2b shows some cooling curves of different quenchants measured by this silver probe, and Fig. 3 shows the heat transfer coefficients calculated from these cooling curves, using the Lumped-Heat-Capacity Method. a) b) Fig. 2. a) Japanese New Silver Probe of 10 mm diameter 30 mm according to JIS K 2242 method B; b) Examples of cooling curves measured by the silver probe [5]

40 Due to the small size of this probe, and the very high heat conductivity of silver, as it can be seen from Fig. 2b, when quenched in oils this probe cools down to 200 C in 15 to 20 seconds, and when quenched in brine, water or polymer-solution of low concentration, in less than 2 seconds. Such probe is well suited for laboratory evaluation of the critical heat flux densities of a quenching fluid in the very beginning of the cooling process. Fig. 3. Examples of calculated heat transfer coefficients from the cooling curves shown in Fig. 2b [5] When quenching real workpieces of bigger size, made of steel, the temperature field across its section is not uniform, and the cooling time to 200 C in the core (depending on the crosssection size and the quenchant's cooling intensity) takes many hundreds seconds, or even several tens of minutes. It is obvious, that the heat transfer coefficient calculated for the small silver probe does not represent the real heat transfer data on the surface of real engineering components, not only because of this vast difference in cooling time, but also because the workshop conditions (fluid flow and direction) greatly differ from those in laboratory tests. 3. TEMPERATURE GRADIENT METHOD FOR HEAT TRANSFER CALCULATIONS WHEN REAL WORKPIECES ARE QUENCHED The method itself, described in detail in [6] is based on the known physical rule that the heat flux at the surface of a body is directly proportional to the temperature gradient at the surface, multiplied by the thermal conductivity of the material of the body being cooled: T q (4) x where: q is the heat flux density (W/m 2 ) that is the quantity of heat transferred through a surface unit per unit time, λ is the thermal conductivity of the body material (W/mK), T/ x is the temperature gradient inside the body at the body surface, perpendicular to it (K/m). When using this method for evaluation of the cooling intensity in workshop conditions, the essential feature is the Liscic/Petrofer probe. It is a cylinder of 50 mm diameter 200 mm length instrumented with three thermocouples placed within the same radius at the half length cross-section. One thermocouple is placed 1 mm below surface, the second one 4.5 mm below surface and the third one at the centre of the cross-section. The probe is made of Inconel 600, the structure of which does not change during heating and cooling, thus there is no latent heat because of structure changes. When the cooling intensity is to be determined, the probe is heated to 850 C until its central thermocouple reaches this value, then transferred quickly to the quenching bath and immersed vertically. The probe is connected to a data acquisition system including a personal computer. The data acquisition card contains three A/D converters and amplifiers with a programme enabling digital recording all three thermocouple signal outputs, and simultaneous drawing three cooling curves in real time, see Fig. 4.

41 Temperature Gradient Method can, of course, be used also for probes of the same design having different diameters, but always the same ratio L/D = 4:1. Fig. 5 shows results of such probes of a) 20 mm Dia. 80 mm and b) 80 mm Dia. 320 mm, quenched in low viscous accelerated oil of 50 C with medium agitation. Fig. 4. Cooling curves measured by the Liscic/Petrofer probe (50 mm Dia. 200 mm) quenched in low viscous accelerated oil of 50 C with medium agitationagi a) b) Fig. 5. Cooling curves measured by the probes of a) 20 mm Dia. 80 mm and b) 80 mm Dia. 320 mm, of the same design as the Liscic/Petrofer probe, quenched in low viscous accelerated oil of 50 C with medium agitation Irrespective of the probe diameter and mass the Temperature Gradient Method exhibits two very important features: a) It displays clearly the dynamic of heat extraction during the whole quenching process. b) It shows the initial heat flux density at the beginning of cooling. ad a) The probe of 80 mm Dia. 320 mm has a mass of 13.6 kg, a ratio surface/volume of only 56 m -1, and a heat capacity of 6045 J/K representing a case of great volume (and heat capacity) and relatively small surface area. The probe of 20 mm Dia. 80 mm has a mass of only 0.2 kg, a ratio surface/volume of 225 m - 1, and a heat capacity of only 94 J/K representing a case of small volume (and heat capacity) and relatively big surface area. The heat capacity of the bigger probe is 64 times bigger than the heat capacity of the smaller probe! Why it is then the complete cooling time to 200 C of the bigger probe (600 seconds see Fig. 5b) only 9.2 times longer than the complete cooling time for the smaller probe (65 seconds see Fig. 5a)? This can be explained by comparing the maxima of relevant temperature gradients: max. ΔT 80 = 415 C and max. ΔT 20 = 114 C, as the mentioned figures show. ad b) Two seconds after immersion the temperature gradient between 1 mm below the surface and the centre of the probe for the small probe was already 29 C, while for the big probe the same temperature gradient was only 8 C, as shown in Fig. 5a, and Fig. 5b respectively.

42 It appears that the smaller probe cools from the beginning faster than the bigger one, but later the temperature gradients within the bigger probe are much bigger (i.e. the heat fluxes are bigger) than within the smaller probe. This is the reason why the cooling time of the bigger probe is only 9.2 times longer than of the smaller probe, although its heat capacity is 64 times bigger. This analysis shows how the Liscic/Petrofer probe, based on the Temperature Gradient Method can precisely describe the dynamic of heat extraction during the whole quenching process. To calculate the heat transfer coefficient based on experiments with the Liscic/Petrofer probe, the 1-D inverse heat conduction method is used. Because of the length to diameter ratio of 4:1 of the probe, the heat transferred through both ends is negligible, and the probe can be considered a long radially symmetric body of a given radius R. The temperature distribution T(r, t) inside the cylinder, for times > 0, depending only on the radial coordinate r from the centre of the cylinder, is determined by the 1-D heat conduction equation: T c div gradt (5) t All physical properties: ρ (density), c (specific heat capacity) and λ (heat conductivity) of the probe's material are temperature dependent, so the whole problem is nonlinear. The initial condition T(r,0) = T o (r) is assumed uniform for 0 < r < R and equal to the initial value measured at the place of the thermocouple. The problem to be solved is to determine the surface heat transfer coefficient α for the boundary condition at r = R: T T Tex (6) x where T ex is the measured external temperature of the quenchant. To determine α, the measured cooling curve at 1 mm below surface at r = r 1 = R-1 mm is used. The inverse problem of computing α is solved by the following numerical procedure: 1. Solve the heat conduction equation (5) within the spatial domain 0 < r < r 1 with the measured T = T mes as a Dirichlet boundary condition at r = r Because r 1 < R, extend the solution towards the boundary from r = r 1 to r = R and 3. Calculate α from equation (6) with measured T ex by using numerical differentiation. Since temperatures are measured at discrete times, they have to be smoothed. This is done by cubic spline least-squares approximation to get sufficiently smooth global approximation over the whole time range. Numerical solution of the heat conduction equation (5) is done by the nonlinear implicit method, with simple iteration per time step, to adjust all physical properties to new temperatures. The solution extension in step 2 is computed by local extrapolation based on low degree polynomial least-squares approximation. The same approximation is also used for the numerical differentiation needed to compute α in step 3. Based on this calculation the heat transfer coefficient is determined as function of surface temperature see Fig. 6a, and as function of time see Fig. 6b.

43 a) b) Fig. 6. a) Heat transfer coefficient for the Liscic/Petrofer probe (50 mm Dia. 200 mm) quenched in low viscous accelerated quenching oil of 50 C with medium agitation, as function of surface temperature; b) Heat transfer coefficient for the Liscic/Petrofer probe quenched in low viscous accelerated quenching oil of 50 C with medium agitation, as function of time 4. CRITICAL HEAT FLUX DENSITIES THEIR INFLUENCE ON DISTORTION OF THE WORKPIECES AT QUENCHING In every quenching process there is an initial heat flux density which depends on the workpiece to be quenched, from the one side, and the critical heat flux densities q cr1 and q cr2 which depend on the quenchant, from the other side. The initial heat flux density depends on the ratio between the volume (heat capacity) and the surface of the body. At the very beginning after immersion, according to equation (4), the heat flux density depends on the temperature gradient at the surface. Bodies having a relatively small volume and a big surface, will have a bigger temperature gradient i.e. a bigger initial heat flux density than bodies having a relatively big volume and small surface, as shown in Fig. 7. Cooling curves measured by the Liscic/Petrofer probe see Fig. 5a and 5b, prove this fact. Fig. 7. Schematic presentation of the temperature gradient at the surface, in the very beginning of cooling a) for a cylinder of 20 mm Dia. 80 mm and b) for a cylinder of 80 mm Dia. 320 mm Critical heat flux densities q cr1 and q cr2 are inherent properties of any vaporizable liquid. The first critical heat flux density q cr1 is the maximum heat flux density that causes film boiling (vapour blanket) at the very beginning of the quenching process, as shown in Fig. 8.

44 Fig. 8. Four modes of cooling at quenching a), critical heat flux densities b), according to [7] The second critical heat flux density q cr2 is the minimum amount of heat energy necessary to support film boiling, this is the point at which the surface of a hot part has cooled enough to allow the collapse of the vapour (end of film boiling), and nucleate boiling begins. There is a relation between q cr1 and q cr2 that is true for all vaporizable liquids: qcr1 5 qcr2 (7) According to [8] upon immersion of a steel part into the quenchant, the initial heat flux density can be: q qcr1; q qcr1; q qcr1 (8) When q q cr1 full film boiling (vapour blanket) will appear. When q q cr1 transition boiling is observed. In case q q cr1 film boiling stage is absent i.e. nucleate boiling starts from the beginning. Each of these three cases will produce different values of the heat transfer coefficient. The first critical heat flux density q cr1 has a great effect on the cooling rate of steel parts and their distortion. It depends on the saturation temperature of the liquid, and the difference between the saturation temperature and the actual temperature of the quenchant. The more resistant a liquid is to boiling, when heat is applied, the higher is the liquid's q cr1. The more resistant a quenchant is to boiling, the more uniformly the part will be quenched (without film boiling) thus yielding less distortion. When water is applied as quenchant the q cr1 value depends on the water flow rate and the water temperature [9]. It can be increased by increasing the agitation rate. Besides, a small amount (e.g. 0.1 %) of chemical additive can increase q cr1 by 2-3 times. To provide for uniform cooling i.e. to eliminate distortion variation, the critical heat flux density q cr1 should be greater than the initial heat flux density. To achieve this, the practical know-how includes the knowledge of the additive, its concentration, and adequate water velocity. Both critical heat flux densities q cr1 and q cr2 can be determined experimentally using a small silver probe as e.g. the Japanese New Silver Probe shown in Fig. 2. Finding the highest q cr1 for a given quenchant will optimize the quench system for all parts quenched in that system, minimizing distortion and maximizing the part properties after the quench. This clearly shows the need to systematically investigate critical heat flux densities for different liquid quenchants in the framework of the proposed database. 5. FACILITIES WHICH ENABLE MEASUREMENT AND RECORDING THE COOLING INTENSITY IN WORKSHOP CONDITIONS When Liscic/Petrofer probe of 50 mm Dia. 200 mm, having a mass of 3.3 kg is used for measurement and recording the cooling intensity of any liquid quenchant (oils, water-based solutions), in workshop conditions, adequate facility is necessary. Besides the required quantity of quenchant it should enable different quenchant's temperatures, and different

45 agitation rates. Fig. 9a shows such facility at the Quenching Research Centre at the Faculty of Mech. Engineering and Naval Architecture, University of Zagreb, Croatia. This facility has a range of working temperatures from 20 C to 80 C, and a flow velocity (agitation rate) from 0 to 1.4 m/s. a) b) Fig. 9. a) Experimental quenching tank of 300 liter capacity for evaluation of the cooling intensity of oils, water, water-based solutions and polymer-solutions b) Experimental saltbath of 1 m 3 capacity for evaluation of the cooling intensity of quenching salts for Martempering and Austempering [10] For isothermal quenching in salt-bath the same centre has a proprietary salt-bath of 1 m 3 salt capacity for Martempering and Austempering processes, see Fig. 9b. By the violent downward flow of liquid salt, a very effective cooling intensity is achieved which is enhanced by automatic addition of small quantities of water. This enables to martemper workpieces of up to 150 mm cross-section, and austemper workpieces of up to 30 mm thickness. The working range of this facility is: temperature 180 to 450 C; agitation rate 0 to 0.6 m/s; water addition 0 to 2 vol. %. 6. CONCLUSIONS Development of new computer aided experimental techniques enable to characterize every liquid quenchant in concrete quenching conditions, in respect of their cooling intensity, more comprehensive and accurate than ever before. The possible consequences of this achievement are twofold: a) Computer modeling of hardness distribution, microstructure, stress and distortion The results of investigations during Phase 1 of the mentioned project, will serve in Phase 2 as input into new advanced 3-D software code for calculation of the heat transfer coefficients at every point of the surface,for real engineering components of complex geometry. This will enable to predict the hardness distribution, microstructure, stresses and distortion at every point of any section of the workpiece. b) Virtual selection of optimal quenchant and quenching conditions Once the Database will contain the mentioned comprehensive information, for sufficient number of different quenchants under specified conditions, virtual computer aided selection of optimal quenchant and quenching conditions, according to specific requirements in every concrete case, will be possible. By the virtual comparison one will gain the following important information:

48 well as multi-processes combined in a single furnace cycle. Technical and technological aspects of the furnace exploitation are presented and operational costs reduction and energy saving are considered. Despite the global economic difficult, the development of vacuum technologies and heat treating equipment continues in applications wherein vacuum is the basis for protective and technological atmosphere. This pertains especially to the applications based on single chamber vacuum furnaces equipped with high pressure gas quench systems (HPGQ). Intensification of gas quench enables heat treatment of not only alloy steels but also steel grades conventionally quenched in oil. There is continuous advancement in vacuum carburizing (LPC FineCarb, PreNitLPC ), which is becoming more and more competitive to traditional carburizing. Furthermore, first applications of vacuum nitriding have appeared (LPN - FineLPN ). Those furnaces are fully automatic, computer controlled and equipped with technical support systems in the form of simulation software for vacuum carburizing treatments (SimCarb, SimHard ) and quenching (G-Quench Pro ). Presently, the HPGQ single chamber vacuum furnaces are capable of handling a number of HT technologies such as: annealing, brazing, sintering, quenching, tempering, carburizing, nitriding, etc. These treatments may be run individually or grouped in a single treatment cycle, e.g. brazing + carburizing + quenching + tempering [1], or quenching + tempering + nitriding, etc. This makes the HPGQ furnace a versatile, flexible and multipurpose piece of heat and thermo-chemical treatment equipment which ensures high quality, repeatability and reliability at a minimal cost, while cutting down on process time and consumption of utilities and maintaining neutrality to the surroundings and the natural environment. SINGLE CHAMBER VACUUM FURNACE HPGQ The sophisticated vacuum furnace HPGQ is a unit featuring an internal quench system based on gas nozzles distributed evenly in the heating chamber around workloads or selectively depending on the shape of the workpieces and workload configuration. The outstanding effectiveness of the nozzle system comes from the fact that the nozzles aim the gas stream directly onto the workload and accelerate to the velocity of km/h. Such an intensive gas stream at high pressure results in very efficient quench and thorough penetration even through densely packed workloads [2, 3]. In the middle of the previous decade the single chamber HPGQ furnaces were only available in the pressure class of bar, which enabled quenching of alloy steels, mainly tool and high speed steels (titanium and molybdenum), for cold and hot working (1.2379, , , , , ), with a limitation imposed by the size of workpieces and workload density. Those furnaces achieved cooling efficiency expressed by heat transfer coefficient α at the level of W/m 2 K. The current standard is class 15 furnaces which feature cooling efficiency of W/m 2 K and thus have a wider application range which Fig.1. Furnace HPGQ 25 bar N2/He type 25.0VPT- 4035/36 manufactured by Seco/Warwick S.A.

49 includes alloy steels for carburizing (16MnCr5, 18CrNiMo7-6), tool steels and HSLA (high strength low alloy 42CrMo4, 40CrNiMo6) for workpieces of small cross-sections [3]. The next border was crossed in 2009 with the appearance of 25 bar single chamber furnaces for nitrogen and helium quench (Fig. 1). These furnaces obtain impressive cooling rates in helium comparable to slow and medium oil quench, at the level of W/m 2 K, which enables quenching of a wide range of typical carburizing steels and HSLA [4] grades and even bearing steels (100Cr6). As far as tool steels are concerned, the cooling efficiency parameter is met even by the standard 10 bar furnaces. Using the example of hot working tool steel H13 (1.2344) and the NADCA [5] heat treatment requirements for dies, a minimum average cooling rate was determined at 28 o C/min in the temperature range of 1030 do 540 o F. The tests following the requirements, done on a 400/400/400 steel block in standard HPGQ furnaces confirmed the effectiveness of the latter in die hardening (Fig. 2). The cooling rates obtained significantly exceeded the limit and, depending on the furnace working area, were respectively: for a 600/600/900 mm furnace approx.. 80 o C/min, 900/800/1200 mm 55 o C/min, and for 1200/1200/ o C/min [4]. Therefore, class 15 bar furnaces may be expected to yield cooling rats higher by approx. 30 % and the 25 bar He ones even twice to three times higher [2]. Cooling rate is of key importance for impact strength and thus for the resistance of tools and dies to thermal fatigue cracking. For example, in the Charpy V-notch test for H13 (1.2344) steel, the cooling rate of 55 o C/min gave the impact strength of 17 J, the rate of 100 o C/min produced approx. 24 J, while a significantly lower cooling rate of 8 o C/min yielded a mere 8 J [6]. In the course of hardening it is equally important to achieve uniform cooling from all sides and to disallow extensive temperature differences between the surface and the core of a workpiece as these may lead to major distortions due to thermal stress and, in extreme cases, cracking or damaging of the tool. It is for such considerations that the HPGQ furnaces are equipped with a system of controlled quench which provides for the adjustment of cooling rate according to one of the thermocouples placed inside a workpiece. Apart from that, Fig.2. Cooling rate test for dies in HPGQ furnace size 600/600/900 mm (Seco/Warwick) Fig.3. The real trend of interrupted quench of the H13 tool it is possible to run cooling based on the temperature difference between the workpiece surface and its core as well as interrupting the surface quench until the core reaches the

50 temperature (interrupted quench, martempering, austempering). The controlled cooling options are provided by workload thermocouples interfaced with fan rotations, which directly influences the cooling gas flow rate. An example of die quench acc. to NADCA with surface and core temperature progress is presented in Fig. 3. A major aspect of enhancing operational properties of the equipment is simulation software which enables prediction of treatment results in the given circumstances. The G-Quench Pro software offered with HPGQ furnaces by Seco/Warwick provides for quench simulations of hot and cold working tool steels. The simulator takes into account a number of quenching parameters such as the type and size of furnace, the type and pressure of quench gas, the steel grade, the geometry of the workpieces and their loading density. Based on the above input, a cooling curve is plotted for a selected point from the surface to the core of the reference part. The outcome of the simulation is a CCT graph with cooling curve and expected hardness (Fig. 4). Furthermore, the software facilitates on-line simulations in real time based on temperature readouts obtained from workload thermocouples directly during the quench. THERMO-CHEMICAL TREATMENT IN HPGQ FURNACES Vacuum carburizing by FineCarb method has been introduced in over 70 industrial applications with single chamber HPGQ furnaces. It is based on a ternary mixture of carburizing gases (acetylene, ethylene, hydrogen) and an adequate manner of treatment, which ensure high efficiency, uniformity and purity [7, 8]. Combined with gas quench it provides an attractive alternative to conventional processes of case hardening carried out in atmosphere furnaces with oil bath [9] (Fig. 5). A further sophisticated stage in the development of FineCarb vacuum carburizing is the currently implemented method of carburizing preceded by nitriding - PreNitLPC. This technology consists in Fig.4. The tool steel hardening simulator G-Quench Pro (Seco/Warwick) Fig.5. Sample workload after carburizing and hardening in vacuum furnace LPC+HPGQ type 15.0VPT-4022/24 (Seco/Warwick) feeding ammonia at the initial phase of treatment i.e. at heating for carburizing. The nitrogen introduced into the surface case in this way aids carburizing by accelerating carbon diffusion and lowering the tendency to build up carbides and, most importantly, by significantly limiting the growth of austenite grain (Fig. 6).

51 Fig.6. The comparison of grain size after the LPC and PreNitLPC treatment at 1000 o C for steel 18CrNiMo7-6 These advantages facilitate a considerable shortening of treatment time through an increase of the carburizing temperature. At the same time the case obtained features a proper microstructure and mechanical properties which are equal to those obtained in conventional treatments at a lower Case Carburising time for 16MnCr5 temperature [11]. For depth comparison, carburizing [mm] 925 o C 950 o C 980 o C 1000 o C 1020 o C 1040 o C with the PreNitLPC 0.5 1h23m 0h57m 0h39m 0h30m 0h24m 0h19m method at the temperature of 1040oC is 4-5 times 1.0 5h30m 3h50m 2h35m 2h00m 1h35m 1h15m shorter than the one h00m 15h10m 10h20m 8h00m 6h10m 4h50m carried out at the conventional temperature of 925 o C (Table 7). 100 % 69 % 47 % 36 % 28 % 22 % Due to their disequilibrous nature, running vacuum carburizing treatments requires a computer assistance. The SimVaC constitutes an integral part of the FineCarb vacuum carburizing technology and of the expert system [12] which focuses on the development of the latter. It facilitates the design of processes of vacuum carburizing and hardening in high pressure gas as well as the analysis and optimization of treatments without the need for real tests which usually are time-consuming and costly. The SimVaC is a sophisticated simulation software consisting of a vacuum carburizing module SimCarb and a hardening module SimHard (Fig. 8). It allows a high precision prediction of the results of real processes based on a process or an outcome simulation. The process simulation presents the effects of a preset process as a carbon profile and a case hardness profile. The outcome simulation Fig.8. The structure of SimVaC simulator vacuum carburizing and hardening (Seco/Warwick prompts the treatment for the input case requirements. The system takes into account the steel grade, the shape and geometry of workpieces, their surface area, carbon concentration in the surface case, the case depth criterion, the carburizing temperature, the time sequence for boost and diffusion. Other factors considered include precooling for hardening, the type and pressure of cooling gas and the size of the furnace. Apart from the carbon and hardness profiles, the simulation yields the demand factor for the mixture of carburizing gases (Fig. 9).

52 Similarly to vacuum carburizing, the HPGQ furnaces may be used for vacuum nitriding (already functioning in the PreNitLPC method). The treatment consists in feeding ammonia to the vacuum furnace chamber at conventional nitriding temperatures. Currently research and tests are being carried out, chiefly on tool steels [13], aimed at mastering the process and working out simulation methods. While it appears obvious that the lengthy nitriding treatments in the HPGQ furnaces will not be justified, an interesting alternative might come from the relatively short, limited to several hours, nitriding of tool steels, applied as complementary to hardening and leading to a very hard and thin case which boosts the functional parameters of the tools. This would be particularly advantageous when the entire heat treatment is done at a single furnace cycle, without opening the furnace door and Fig.9. The outcome of SimVaC simulation of vacuum carburizing (LPC) and gas hardening (HPGQ) shown as a hardness profile based on carbon profile. During the 2 h treatment at the temperature of 1040 o C the case obtained was 1.40 mm for steel 20MnCr5 transferring the workload, by going through a sequence of: hardening, multiple tempering and final nitriding. Further advantages of the so conducted treatment are: excluding the chemical activation of surfaces before nitriding and obtaining a rapid and uniform increase of the nitrided case. This is due to heating in vacuum, which has strong reduction properties and cleans and activates the surfaces of the workpieces. To confirm the above, a complete heat treatment of steel tools was run in a HPGQ furnace type 15.0VPT-4022/24. Fig.10. The trend of complex heat treatment process of tool steel H11 Fig.11. The hardness profile and microstructure obtained after complex treatment for steel Austenitization was effected at the temperature of 1030oC, followed by hardening in 12 bar nitrogen, then twice tempered at 570oC/2 h and finally nitrided at the temperature of 540oC for 4 h (Fig. 10). The treatment resulted in a uniformly nitrided diffusion case of approx. 0,14 mm, surface hardness of approx. 900 HV and core hardness of 500 HV, respectively (Fig. 11).

53 ENERGY EFFICIENCY A reduction in energy requirements for the HPGQ furnace occurs at a few areas simultaneously. Among the basic areas are technical solutions which reduce thermal losses in the heating chamber and the application of electric receivers of increased energy efficiency. Very important is the optimization of the process, mainly its length. For that purpose a temperature monitoring in the workload is used as well as processes of high temperature carburizing (PreNitLPC ). As far as electrical power supply is concerned, it is essential to ensure a stable power demand not exceeding the maximum level while maintaining the highest possible power factor (P/S). The HPGQ vacuum furnaces are equipped with a power management system which ensures: - Power demand limitation depending on temporary requirements. - A substantial improvement of the power factor P/S in the heating and cooling phases. - Start-up of fan motor without exceeding the rated currents (elimination of the starting current peak in the motor). - Increased efficiency of the blower motor. The HPGQ vacuum furnace comprises two main systems which use up most of the electrical energy (heating and cooling), each of which is conventionally equipped with individual power control systems. For heating, there are SCR controllers or a transducer to control the power of the resistance heating elements. For cooling, there are a soft-start or an inverter to control the blower motor. Since the furnace operating sequence does not provide for simultaneous heating and cooling, only one of the control systems may be activated while the other one is switched off. This dependency led to a search for a single system capable of alternate control of heating power or cooling intensity in the furnace. After theoretical analysis and testing an appropriate device was found an inverter which, apart from controlling the work of an induction motor, may control the power of the resistance heating elements supplied through a transformer. What is more, such application enhances the operating features of the furnace and decidedly reduces its power consumption by increasing the power factor PF=P/S. Fig.12. The demand for active, passive and apparent electrical power during a hardening process depending on heating control mode: conventional thyristor (SCR), and inverter (INV)

54 Tab.13. The energy comparison for reference treatment for thyristor (SCR) and inverter (INV) controllers Energy consumption SCR INV Safe Pt [kwh] Qt [kvarh] % St [kvah] % PF av % Fig. 12 presents a comparison of electrical power requirement for active, passive and apparent power during the hardening treatment as exists between the traditional SCR controller and an inverter. In the given case the inverter control reduced energy consumption by 42 % and improved the power factor by 27 % (Table 13). This method of heating and cooling with the aid of an inverter is protected with a patent [14] and has been successfully used in a few dozen HPGQ furnaces all over the world. SUMMARY The single chamber vacuum furnaces HPGQ by Seco/Warwick provide a versatile and efficient tool for heat and thermo-chemical treatment. They are sophisticated devices which comply with the toughest quality, economic and environmental standards thanks to the following advantages: - Multiple increase in cooling rate in the class of bar N2/He furnaces enables heat treatment of steels conventionally hardened in oil. - The exclusion of quench oil eliminates washing and utilization of washing means and the oil, which reduces the HT cost and makes the technology environmentally friendly. - The single chamber furnace within which the workload is not moved enables application of workload thermocouples and complete temperature monitoring inside the workpieces, which in turn permits process optimization while simultaneously meeting the stricter requirements (the aviation industry). - The function of isothermal and controlled cooling enables control of cooling rate and temperature distribution in the workload, thus reducing potential deformations and the risk of thermal fatigue cracking. - The advanced high temperature carburizing acc. to the PreNitLPC method permits a multiple shortening of treatment time, thus minimizing the costs. - The possibility of vacuum carburizing and nitriding combined in a single cycle with hardening and tempering broadens the technological potential by including the multistage processes in one cycle. - The simulation software for hardening and carburizing enable prediction of treatment results with high accuracy and eliminate the need for tests done on the treated workpieces. - The electrical power management system facilitates economical and optimal use of energy. REFERENCES [1] KORECKI M., ADAMEK A., Flexibility and Versatility of Heat Treatment under Vacuum. Furnace International, Part 1 March/April 2005, s. 6-8, Part 2 May/June 2005, s. 8-12

63 reproducibility benefits. This technology is introduced as an economical complementary solution in vacuum heat treatment when the same high quality material is required. Other information on B.M.I. Fours Industriels

65 efficiency of industrial furnaces. Investments provide a double benefit: for the operating costs and for the environment. The most environmental friendly and safest kilowatt hour is the one that is not consumed (German environment secretary Sigmar Gabriel). Nowadays the topic energy efficiency spotlights not only due to the check of possible saving potentials during the financial and economy crisis but also through new laws. A new Energy using Products -regulation (EuP regulation) currently passes the authorization procedure of the European Union. This will then have to be observed by the manufacturers of thermoprocess furnaces besides the already existing Eco-Design-Regulation 2005/32 EC in future. 1.1 Heat energy processes during the heating cycle within the vacuum furnace At any kind of heat treatment process not only the heat treated parts with the corresponding loading media but also the heating elements and/or the complete hot zone is heated to the necessary process temperature. A high extend of energy efficiency has thus to be pushed during the heating mechanism. Also at heating processes in graphite insulated vacuum hardening furnaces (pic. 1, pic. 2) a high extend of the fed energy is lead into the dead masses of the hot zone. The bigger the hot zone s volume and the load the more energy has to be fed in for the heating. The aim is of course to run as much parts as possible within one cycle. Thus the volume s reduction is restricted only to the design features of the hot zone and where applicable to the loading media. Pic. 1 SCHMETZ one-chamber vacuum furnace Pic. 2 Hot zone with graphite insulation Graphite material is mainly used for the insulation of hot zones of vacuum furnaces due to its high temperature and form resistance. All parts of the heating are optimised regarding their functioning and weight. The heating rods and bridges weigh as less as possible without having any risks regarding their stability and life time. During the heating and temperature soaking cycle (pic. 3) the fed electrical heat energy has to compensate dead losses. These dead losses of the hot zone are reduced through an insulation structure that is as tight as possible. Usually a 40 mm thick graphite felt plate is installed as basis. The characteristics of each graphite felt material play a decisive role. Many different insulation characteristics can be noted in the several material qualities offered at the market which laymen are not able to distinguish. Thus the exclusively use of OEM-quality is absolutely necessary.

66 Pic. 3 Vacuum furnace: Convection heating process By using increased insulation better insulation values are achieved. An increase of the hot zone s insulation from the usual 40 mm (for example SIGRATHERM ) to 60 mm reduces the dead losses in temperature soaking cycles by approx. 15 %. 1.2 Heat energy processes during the cooling cycle within the vacuum furnace Also at the cooling cycle the complete hot zone is cooled down besides the load (pic. 4). The necessary energy expenditure should as well be reduced as far as possible. This means the actively to be cooled dead masses have to be minimized. Pic. 4 Vacuum furnace: Cooling process For the cooling of the load not only the cooling gas pressure and speed are important but also the gas flow rate and gas distribution throughout the load space. To guarantee a uniform and distortion-less quenching the gas flow is distributed via special gas guiding devices throughout the complete load space. The gas guiding devices integrated in the hot zone gas distribution plates made of hard graphite (pic. 5) have to guarantee an optimum flow ratio on the one hand and on the other hand have to meet the requirements of a minor dead weight at simultaneous high life times.

67 Pic. 5 Standard gas guiding system made of hard graphite At the gas guiding devices new developed insulated nozzle plates (pic. 6) can be installed instead of the so far used distribution plates. The new developed nozzle plates have the advantage of a considerably low net weight. This means for example a mass reduction of 26 kg at a standard vacuum furnace with the dimensions of useful space 600 x 900 x 600 (w x l x h) with one gas inlet and gas outlet opening each. Here the energy need of the dead mass is reduced at the heating as well as at the cooling process. Pic. 6 Optimised gas guiding systems with nozzle plate 2.0 ENERGY SAVING SYSTEM *ESS* To achieve a high degree of energy efficiency at the heating as well as the cooling the SCHMETZ system *ess* (energy saving system) combines the plant-specific improvements of an increased hot zone insulation and the weight-optimised nozzle plates as gas guiding devices. Besides the process temperature the current consumption in the vacuum heat treatment process generally also depends on indicators like for example heating ramps, soaking times, cooling gas pressures, revolution number of the motor, a.s.o. A heating to a high hardening temperature (for example high-speed steel) and a rough overpressure gas quenching have a relatively high current consumption. A heating process to a low tempering temperature (for

68 example stainless steel) or a slow cooling (for example at vacuum brazing processes) however needs considerably less electrical energy. The degree of the current consumption reduction by means of the energy saving system *ess* thus depends on the application. In addition it varies in the single cycles of a heat treatment process. By means of the mentioned effects at the heating and cooling, vacuum furnaces with the innovative hot zone design SCHMETZ system *ess* achieve savings of about % of the normal current consumption. Simultaneously shorter heating times are realized. A hardening shop of a known group of companies operates among others two vacuum hardening furnaces with the identical dimensions of useful space 900 x 1200 x 700 mm (w x l x h). The older furnace has a usual graphite insulation thickness and standard gas distribution plates, the new furnace is equipped with the energy saving system *ess*. Identical hardening and tempering loads with 1500 kg tool steel were heat treated in both furnaces. At this very good load capacity a current saving of 362 KWh 14,6 % could be detected with the new system for the complete hardening and three-times tempering. Simultaneously the process time was shortened with about 4 hours by 10,7 % (pic. 7). Pic. 7 Heat treatment cycle load kg, hardening and 3x tempering in a vacuum furnace with standard design and a vacuum furnace with the energy saving system *ess* 3.0 ADDITIONAL CAPACITY At heat treatment processes of different kinds and in different sized furnaces a considerable process time reduction could be determined. In the above mentioned 36h-hardening and tempering process the process time is reduced by 10,7 % to about 32 hours. The operator of these two furnaces has an average effective furnace utilization of about 5800 h per year. This means that in this standard furnace for this process with 36 h operating time approx. 161 loads per year would be possible. At a shortened process with an operating time of 32 h with the system *ess* this can be increased to approx. 181 loads per year.

69 4.0 MAXIMISING OF THE COOLING SPEED AT THE PART AND OPTIMIZING THE QUENCHING HOMOGENEITY The weight optimised gas guiding devices with inflow nozzles (nozzle plates) achieve a higher quenching speed at the part. Comparative measurements were carried out in corresponding furnaces with the dimension of useful space 600 x 900 x 600 mm (w x l x h) and a 340 kg (gross) bolt load. Thermocouple measurements in reference bolts of different diameters prove a cooling speed increase of approx. 10 %. In addition a higher quenching homogeneity could be proven in analysis at the Berner Fachhochschule HFT Biel, Switzerland, also at a smaller vacuum furnace with the dimensions of useful space 400 x 600 x 400 mm (w x l x h) with the new nozzle plate compared to the standard gas distribution plates. 5.0 PROFITABILITY AND DEPRECIATION The European industrial current prices vary between the local suppliers and highly depend on the ordered quantity. For example: one bigger contract heat treatment shop has a typical current price (incl. a power ratio) of 0,12 per kwh and a smaller in-house heat treatment department has to pay 0,18 per kwh. The profitability of the *ess* system is analysed at an example of a vacuum hardening furnace with the standard dimensions of useful space 600 x 900 x 600 mm (w x l x h). At the hardening process at this furnace size with standard features an average current consumption of about 70 kwh per hour is assumed. For a vacuum hardening furnace of this size and a good capacity utilization of 7000 furnace operating hours per year a current consumption of 490,000 kwh can be assumed. At an expected current price of 0,15 per kwh (incl. power ratio of the supplier) the current costs for the hardening operation of this standard furnace would be 73, per year. Thus an average current consumption reduction of 15 % saves 11, per year here. The equipment with increased hot zone insulation and nozzle plates for this furnace size additionally costs about 25, The depreciation time would thus be < 2,5 years. The life time of the graphite part of the hot zone can be influenced by several factors (for example burn-off through oxygen break-in, mechanical wear and tear caused by not cleaned parts, bring-in of cuttings, damage during loading). But providing appropriate operating and thorough furnace care an average hot zone life time of eight years can be reached. The total cost saving for the complete hot zone life time would be in this example by means of the reduced current need 8 x 11, = 88, Considering the additional investment costs this means a profit of 63, ( 88, ,000.00). With this the occurring costs for the hot zone replacement can be covered for example after eight years of operation. To complete the efficiency consideration the already mentioned increased furnace reliability has of course to be considered. Thus a higher load volume can be treated with the invested furnace. 6.0 RETROFIT ENERGY EFFICIENCY The SCHMETZ system *ess* was installed and commissioned in new vacuum furnaces with different standard dimensions of useful space internationally. But the new concept with increased hot zone insulation and gas distribution with nozzle plates can also be retrofitted in

70 almost all older SCHMETZ one-chamber vacuum furnaces. Especially in the course of a hot zone replacement that has to be carried out anyway in OEM quality a retrofit pays of environmentally and financially very fast. CONCLUSION Vacuum furnaces with the innovative hot zone design SCHMETZ system *ess* achieve less current consumptions % of the usual current consumption and thus direct operating costs can be saved. Process time reductions of 10 % can simultaneously be achieved and thus a considerable increase of capacity can be realised. In addition weight optimised gas guiding devices with inlet nozzles (nozzle plates) achieve a possibly faster and even more homogeneous quenching speed at the part. Comparative measurements show a cooling speed increase by approx. 10 %. Furnace operators as well as heat treatment customers can financially participate in the energy efficient furnace technology. With regards to the environment the following is valid: The most environmental friendly and safest kilowatt hour is the one that is not consumed.

96 VLIV PULZNÍ PLAZMOVÉ NITRIDACE PulsPlasma NA NÁKLADY A PROSTŘEDKY TEPELNÉHO ZPRACOVÁNÍ POVRCHOVÉ VRSTVY U PŘEVODOVEK A NÁSTROJŮ COST- AND RESOURCE EFFECTIVE SURFACE LAYER HEAT TREATMENT IN GEAR AND TOOL INDUSTRY BY PulsPlasma -NITRIDING Reinar Grün a, Dietmar Voigtländer b a PlaTeG GmbH, Siegen, Germany, Fon: , b PlaTeG GmbH, Siegen, Germany, Fon: , ABSTRACT Pulzní plazmová nitridace PulsPlasma převodovek a nástrojů se čím dál tím víc pouņívá jako alternativa ke standardním procesům tepelného zpracování pro zvýńení tvrdosti povrchové vrstvy, pro zlepńení otěruvzdornosti a odolnosti proti korozi. Díky specifickým podmínkám procesu tím dochází ke zvýńení ņivotnosti součástí oproti tepelnému zpracování vysokoteplotní cementací nebo standardní nitridací plynným amoniakem. Pulzní plazmová nitridace PulsPlasma je bezodpadová technologie, která ńetří energii a dalńí prostředky. Celkové náklady na výrobek je moņné významně sníņit. Díky vysoce zdokonalenému zařízení pro technologii PulsPlasma je moņné jednotně zpracovávat jak mnoho malých dílů v jedné vsázce, tak větńí nástroje a díly převodovek pro větrné elektrárny. The PulsPlasma nitriding of gears and tools is used more and more as an alternative to standard heat treatment processes for surface layer hardening for the improvement of wear and corrosion protection. By this the lifetime will be longer due to the specific process conditions against the case hardening by high temperature carburizing or by standard Ammonia nitriding (gas nitriding). The PulsPlasma nitriding is a pollution free technology saving energy and other resources. The total manufacturing costs of workpieces can be reduced significantly. Highly developed concepts of plants and the use of PulsPlasma technology a uniform treatment of many small parts in one workload is possible as well as the nitriding of large tools or gear parts for wind power plants. 1. INTRODUCTION For the improvement of wear resistance of tools and gear components of steel the functional surfaces are case hardened usually. Depending on the specific application and the component of expected load the designing engineer defines both the material, as well as the process characteristics like surface hardness and case depth. That means, for example for wind energy gears subject to high stress, these gear wheels must be carburised at temperatures of more than 900 C longer than 90 hours before the hardening in order to realize case depths between 1 and 2 mm. The case hardening leads to structural

97 changes in the treated material and therefore to measure and shape changes. The handled components must be kept on in an additional heat treatment step in order to reduce the internal stress. For the adjustment of the required surface quality and the final dimensions an extensive mechanical processing of the parts after the heat treatment is needed. An alternative to the case hardening usable surface layer heat treatmentprocedure is the nitriding. It is a thermo chemical diffusion method for the enrichment of the workpieces surface zone with Nitrogen. This accepts in this case chemical compounds with the base material and his alloy components. As a result of the nitriding treatmentarises in the peripheral zone a nitriding layer with an external area, the so called white layer or compound layer (CL), and in the direction to the core subsequent diffusion zone (DZ). Accompanied through a nitriding process is the formation of a hardening zone in the border area and the formation of internal compressive stresses through the procedure characteristic. This will allow to influence the wear behaviour, the corrosion resistance and the endurance of a workpiece without changing the materialdependent mechanical properties of the component. Indeed, a major advantage of nitriding in contrast to the case hardening is, that the necessary heating for diffusion of nitrogen can be made below the tempering temperature of heat-treated- and tool steel. Microstructure transformations and the consequent loss of dimensional stability and strength are avoided during nitriding. Nitrided parts are immediately usable and require, with a few exceptions, no post machining. 2. NITRIDING PROCEDURE OVERVIEW Nitriding processes are often referred to the physical state of aggregation for the nitrogen donor starting compounds. - liquid: Salt bath for nitrocarburising - gaseous: Gas nitriding, for ammonia gas nitriding and nitrocarburising - ionised gas: Plasma nitriding, Plasma nitrocarburising The aforementioned nitriding procedures have their pros and cons, which before deciding in favour of a possible nitriding alternative to the case hardening in terms of the required component properties and accessible nitriding characteristics must be balanced. Nitriding Method Salt Bath Nitriding Gas Nitriding Plasma Nitriding Nitriding Medium Cyanide/ Cyanate NH 3 NH 3 + CO 2 N 2 + H 2 N 2 + H 2 + CH 4 Treatment Temperature ( C) Duration of Treatment (h) Result (480) ,2 3 Carbonitrides , ,2 6 Nitrides Carbonitrides Nitrides Carbonitrides Table 1: Comparison of various nitriding methods

98 2.1 Salt bath nitrocarburising The salt bath nitrocarburising is a very flexible nitriding procedure due to the short processing times at higher temperatures. It is usefully usable on all those applications, at which in the first place it depends on the rise of the wear resistance and/or corrosion resistance of the component surface. More or less serious procedure disadvantages restrict, however, the applicability in particular for the nitriding treatment of large components: - High washing effort after the nitriding-treatment, - High regeneration and disposal costs for salt and washing lye, - High energy costs for the operation of a bath, therefore restricted bath size, - Treatment temperature strongly restricted, - Partial nitriding is difficult, 2.2 Gas nitriding and gas nitrocarburising Gas nitriding and gas nitrocarburising are universally applicable nitriding procedures, that have experienced an plant specific and control technical further development in the last 10 years. Both process variants are very well alternatively usable to the case hardening. In particular during the treatment of big tools and gear wheels are to be expected due to the clearly smaller treatment temperatures as well as the immediate suitability for further processing of the nitrided economic advantages components opposite to the case hardening. In spite of the achieved high technological level of these procedures are resulting from the treatment needful raw gases Ammonia and Carbon dioxide and some procedural basic conditions disadvantages of process engineering that limit the applicability of the technical, economic and environmental point of view in some cases: - High gas consumption (m 3 /h) - Use of flammable gases, i.e. special measures are required to ensure a save kiln operation - No possibility to remove the natural passive layers and/or. coming passive layers from the manufacture of parts during the nitriding, - Nitriding of rust- and acid-resistant steels not possible, - High expenditure for covering areas that are not to be nitrided and after nitriding for removing the cover paste 2.3 PulsPlasma -Nitriding First applications of the plasma nitriding go back until the thirties and forties of the last century. Later, in the sixties/seventies, the procedure was developed up to the industries requirements. First direct current cold wall installations for plasma nitriding were used in production. Another development was the thrust of the plasma nitriding mid eighties with the introduction of the so-called pulse technique. This is done by means of plasma excitation in a pulsed DC area. The emergence of arcs is determined by the constant interruption of the voltage flux is avoided. And last, but not least, the pulsed DC voltage allows the necessary heating up to the treating temperature and decoupling from supporting plasma power. That necessary equipment for the cooling of the vessel wall to transfer the excess heat energy (cold wall installation) can be omitted. Hot wall installations with separate heating of the vessel wall are now state of the art

99 in plasma nitriding. With the classical salt-bath nitrocarburising with cyanide containing salts and the gas nitriding with ammonia the dissociation effects of the Nitrogen-giving donor media and the formation of the nitride in a thermo chemically activated process under atmosphere pressure conditions and/or in the low pressure area. The energy needed for the activation of the process and for the formation of nitrides is provided on thermal ways. For the maintaining of the nitriding a certain minimum temperature is necessary, below that temperature the nitriding process cannot work or it will be uneconomically slow. Process temperatures are mentioned in table 1. In Contrast to this the energy of a high-current glow discharge is used by the PulsePlasma - Nitriding in order to activate the reactions necessary for the surface layer, the dissociation of the nitrogen molecule into his atoms, ions and electrons. The parts to be nitrided (charge/workload) are loaded into a external heated vacuum chamber. (hot wall system). After the evacuating of the chamber on working pressure (50 to 400 Pa) a pulsating voltage of several hundreds of volts is applied between the charge (cathode) and the chamber wall (anode), so that the process gas in the chamber is ionized and becomes electrically conducting by this. Depending on the applied voltage between the components to be treated and the chamber wall the ignition of the glow-discharge will be done. Such a glow has a characteristic light which depends on pressure-, temperature- and gas-mixture. Fig 1: Glow discharge around a tool during PulsPlasma -nitriding The reactive Nitrogen ions in the gas mixture can form a chemical compound with the iron atoms of the steel to be nitrided. In addition to this nitrogen atoms will diffuse into the surface layer of the steel. This diffusion is depending on temperature and treatment time. The diagram of an installation for PulsPlasma -Nitriding is shown in picture 2.

100 Fig. 2: Diagram of a PulsPlasma -nitriding plant To For the PulsPlasma -Nitriding and/or -Nitrocarburising Nitrogen /Hydrogen gas-mixtures and gases with Carbon additions, like Methane, are used. By the nitriding the formation of a compound layer Fe x N y (iron nitride) at the surface of the components will be done. Depending on treatment time and treatment temperature nitrogen atoms diffuse into the surface layer of component and will form the diffusion zone. This can be either atomic nitrogen, dissolved in the iron lattice, as well as in the form of included nitride precipitations. PulsPlasma -nitrided layers have basically the same structure like nitriding layers, which are produced by other nitriding processes. The thickness of compound layer will depend on material quality and process parameters in the range between 1 to 20 µm. The thickness of the diffusion zone, characterized by the nitriding case depth can reach up to 0.6 mm under standard nitriding conditions (picture 3). Fig. 3: Surface layer of a nitrided steel component The nitriding of components with a case depth higher than 0,6 mm, for example for gears is possible by using the right material with adapted process parameters for plasma nitriding. subject to high stress, is in the case of choice of suitable materials possible.

101 3. PROCESS ADVANTAGES PulsPlasma -NITRIDING 3.1 Temperature distribution The use of a chamber with an electrically heated wall has a considerable influence on the temperature distribution within a charge next to energy saving-effects due to the wall insulation. In order to avoid a significant temperature rising in the centre of a charge of a cold wall chamber, people refrain in many cases from using this room. Therefore the charge is arranged in a ring-shaped way in a cylindrical chamber. With hot wall installations a nitriding treatment can be carried out through the reduced energy supply about the pulsating plasma for a completely filled charge without the risk of the overheating at certain places. Both the nitriding of charges closely packed and also the treatment of big components is realizable due to the good temperature uniformity in PulsPlasma -Nitriding installations (picture 4). Fig. 4: Various Gear wheel loads Bell type PulsPlasma -Nitriding installations are used mainly due to handling advantages compared to pit type installations (picture 5). Fig. 5: PulsPlasma -nitriding plant Ø 2000x3000 Tandem design The charge is built up on an available base directly or with the aid of an workload fixing.

102 Such a work load fixing can be filled and prepared outside of the chamber and moved into it by a crane. In the case of especially big, heavy tools or gear components it can be reasonable to use another furnace principle and to carry out the installation as a chamber furnace with a front door. Such a PulsPlasma -Nitriding furnace for the treatment of sidewall tools with unit weights of up to 40 t is shown in picture 6. Fig. 6: PulsPlasma -nitriding plant Ø 4300x10500 Chamber design With such a installation a truck-type charge wagon is loaded by means of crane with the tools to be nitrided and they are moved into the chamber together with this. In this way big and heavy components are easy to handle and will be very good for the nitriding process inside the chamber. 3.2 Process gas consumption PulsPlasma -Nitriding Nitrogen-Hydrogen-Methane gas mixtures are used, depending on the application and the results required. During the nitriding no harmful reaction products will be formed so that the process exhaust gases can be given to the environment. Due to the plasma activated nitriding process and the fact of working at low pressures the process gas consumption is very low. An installation with the chamber size of Ø1200x2000 mm height requires an average of 180 l/h process gas mixture. A gas nitriding installation with similar chamber size would have to be supplied with 6000 to l/h ammonia and carbon dioxidecontaining process gas mixture. For classical case hardening processes the required gas volume will be as high. Due to these facts, the ammonia-gas nitriding and case hardening are producing great amounts of waste gases inflammable and harmful which have to be disposed under expenditure of additional energy, for example by post-combustion. 3.3 Variable treatment temperatures Due to the additional plasma activation of the nitriding process and the dosing of the plasma power by the pulsated mod it is possible to carry out PulsPlasma -treatments in a wide temperature range between 350 C and 600 C. Parts with a risk of distortion can be nitrided under optimal low temperature conditions. The core hardness and themechanical strength of

103 the nitrided components is maintained because nitriding temperatures can be used which are below the tempering temperature of the base. After the nitriding no further heat treatment is necessary. PulsPlasma -nitrided components can be used immediately. Steels with higher Chromium content can be nitrided by salt bath only under loss of the corrosion resistance and by gas nitriding only with an additional depassivation procedure before nitriding. By PulsPlasma -Nitriding such a treatment can be done directly. Before the nitriding treatment the required depassivation of the surface will be done by the ionic bombardment of the surface in a process-specific way. By choosing nitriding temperatures below 450 C and with exact controlling of the gas mixture the nitriding requirements are adjustable and controllable. This allows to form hard, wear-resistant nitriding layers at the surface of the components without reducing the corrosion resistance of the material. 3.4 Treatment of sintered steel Components made of sintered steel can be treated by case hardening. Salt-bath nitrocarburising and gas nitriding/-nitrocarburising are possible, but due to the porosity and the process conditions a hardening inside of the sinter material is done also. The complete component could be get brittle in result. During the treatment in the plasma only the outside surface, surrounded by the glowdischarge, is treated. Due to the low pressure (vacuum) and the low gas flow during the plasma nitriding the risk of an over nitriding inside the material does not exist. Any calibrating oil or similar inside the porosity of components have to be removed before a nitriding-treatment will be done. A subsequent treatment and/or cleaning of the parts after the plasma nitriding is not necessary. As an example gear parts made of sintered steel are shown in picture 7 after a PulsPlasma -Nitriding treatment. Fig. 7: PulsPlasma -nitrided sintered components 3.5 Partial treatment No other surface layer heat treatment procedure offers so simple possibilities of the partial treatment as the PulsPlasma -Nitriding. Areas that may not to be nitrided can be masked mechanically by steel plates, tubes or bolts. A painting with specific pastes can be used also,

104 but have to be removed after the treatment. The surface quality of the covered areas remains unaffected., in which only the threads at the ends had to be treated for the reduction of the galling. This could be achieved by a tube covering of the piston-rod area that was also the condition for the surface rawness of this face should not influenced. Fig.: 8 PulsPlasma -nitrided Filter rod threads 3.6 Process combinations Due to related procedures and almost identical systems engineering it is possible to combine different surface treatments in a PulsPlasma -unit in one process. For the further improvement of the corrosion resistance of nitrided components a PulsPlasma -Oxidation treatment can be connected to a PulsPlasma -Nitridingt, just by changing some process parameters and process gases. This oxidation treatment will form a 1 to 3 µm thick Fe 3 O 4 - layer on the compound layer. Depending on the steel quality and the used nitriding process a corrosion resistance of up to 200 hours can be achieved in the salt spray test according to DIN or SAE. A further advantage of the subsequent oxidation treatment is to improve the sliding properties of the treated surface, so that under certain circumstances, the lubricant used for such friction pairs treatment can be reduced in such a way. A further alternative is a so called duplex coating, the combination of PulsPlasma -Nitriding with a Plasma CVD-hard coating or DLC deposition (Diamond like carbon), which are only few µm thin. By the first nitriding treatment a good support of the extremely hard, wearresistant Plasma-CVD coatings will be reached. A further result of such a treatment is the increase of durability for tools and machine parts. An Picture 9 shows as one example an extrusion tool for plastic production after a duplex treatment of PulsPlasma -Nitriding and PulsPlasma -CVD-DLC-coating.

106 An aspect wider, not to be underestimated is added: The economical one. An industrial example is shown as follows. Sometimes it is absolutely reasonable to rearrange the part production in order to be able to reduce the energy and costs-intense case hardening of gear components for the benefit of a PulsPlasma -Nitriding. It has to be considered, that surface layer properties as surface hardness, wear resistance, are similar or even better compared to the case hardening. After case hardening a lot of parts will have to be reworked due distortion by use of high temperature and quenching for hardening. In contrast to that the plasma nitriding can be done at lower temperatures with minimized distortion. Strength demands, which in combination with the selected thermal treatment methods on the wear and lifetime behaviour of a gear, may require by selecting an appropriate base material also can be done by means of nitriding. In the reported example [1] should replaced the case hardening of gear wheels for printing machines made from case-hardening steel 15 CrNi 6 E by PulsPlasma -nitriding. First of all a suitable material was determined for this purpose on calculative way and through practice tests. Picture 11 gives an overview of the calculated material characteristics. Fig. 11: Strength calculation of gear wheels from various materials As a result of the change of different production steps savings of expenses up to 30 % could be realized next to a better service performance of the gears in the component manufacturing.

107 Fig. 12: Manufacturing costs for various heat treatment methods 5. SUMMARY The PulsPlasma -nitriding of components for the improvement of the wear and corrosion behaviour and to extend the life time is described and explained. Due to some further procedural advantages this process is increasingly used for surface heat treatment. In contrast to the case hardening this nitriding process makes a resource saving heat treatment possible and leads very often to the lower manufacturing and operating costs as a whole. Technically sophisticated installation concepts and the use of PulsPlasma -Technology allows both, a uniform treatment of large quantities of smaller components in a charge, as well as the nitriding of large tools or gear wheels. Through process combination with a post oxidation treatment or a plasma-assisted hard coating and/or -DLC deposition in one installation, component properties keep on being able to be optimised. 6. LITERATURE [1] Urban Spatz, strength test of hardened and plasma nitrided gears wheels for printing machines, final year project, University Augsburg 1995

109 bandings, clusters or other inhomogeneities [1-3]. These facts are displayed in excellent mechanical properties, particularly in toughness and fracture toughness [4]. Plasma nitriding is a widely used surface treatment technique bringing an increased hardness, wear resistance, reduced friction coefficient and better corrosion resistance. These properties can be attributed to the presence of fine nitrides, which are formed via the interaction of nitrogen atoms with iron and alloying elements [5-11]. However, hard surface region is sensitive to the initiation and the propagation of brittle fracture since it can be expected that it s fracture toughness is very low. Current paper deals with the evaluation of toughness of no-nitrided and nitrided Vanadis 6 ledeburitic tool steel and brigs some new remarks and recommendations to the practice, to end-users of nitrided tools. 2. EXPERIMENTAL Specimens with various cross section, with a 100 mm in length, made from the ledeburitic VANADIS 6 steel (2.1 %C, 7 %Cr, 6%V, Fe bal.) were heat treated (austenitized, nitrogen gas quenched and double tempered) in a vacuum furnace to a resulting hardness given in Table 1. Subsequently, the specimens were plasma nitrided in the RUBIG Micropuls - plasma equipment. Various combinations of the temperature and dwell time were used, see Table 1. In this Table, important parameters of nitrided region are recorded, also. Table 1 Surface content of interstitials, nitrogen diffusion depth and carbon redistribution vs. nitriding parameters. Nitriding Nitrogen surface content (wt.%) Carbon surface content (wt.%) Nitrogen diffusion depth ( m) 470 o C/30 min o C/120 min o C/60 min o C/120 min For microstructural examination (including the fractography) the light microscopy, the scanning electron microscopy and the high resolution transmission electron microscopy (HRTEM) were applied. Samples for light microscopy were standardly prepared, e.g. ground and polished with diamond paste, the last step with the 1 m diamond paste and etched with Nital-reagent and Murakami-reagent, respectively. The specimens for the scanning electron microscopy were prepared by deep Nital-reagent etching. For the HRTEM, thin foils were cut from bulk samples and prepared by Ion Slicer. The depth profiles of carbon and nitrogen were established with the WDX - microprobe analyser CAMEBAX-MICRO. Five measurements were made on each specimen and the mean values and standard deviations were calculated. Hardness measurements using a Vickers s method were performed using a Zwick hardness tester, applying a load of 10 kg for surface hardness and 50 g for hardness depth profiles determination. The depth profiles were measured on carefully prepared specimen cross-sections at an inclination angle with respect to the surface (bevelled specimens) of 12 o, to improve the depth resolution. The distance between the adjacent indents was 20 m. The toughness has been investigated using the three point bending test method. The distance between supports was 80 mm. Specimens were loaded in the central region, at a loading rate of 1 mm/min, up to the moment of the fracture. Five specimens were tested at each combination of nitriding parameters.

110 3. RESULTS AND THEIR DISCUSSION Figures 1 a and 1 b show that the microstructure of the substrate material consists of the matrix with fine, uniformly distributed carbide particles of regular, spherical shape and with a size of several microns. The carbides are of two basic types. The first ones are larger, white regular particles of the M 7 C 3 -phase. The second type is the MC phase, visible as smaller grey spherical formations. The matrix consists of needle like tempered martensite, Fig. 1b. As reported previously [12], tempered martensite is of two basic types: the dislocation type with a high density of dislocations and the twinned martensite. No proeutectoid phases were identified at the grain boundaries. M 7 C 3 a 10 m b 3 m Figure 1. Light (a) and SEM (b) micrographs of the material after heat treatment 2 10 m 3 10 m 4 10 m 5 10 m Figures 2-5. Light micrographs showing the microstructure of nitrided regions developed at: (2) 470 oc/30 min., (3) 470 oc/120 min., (4) 500 oc/60 min., (5) 530 oc/120 min.

111 Light micrographs in Figs 2 5 show the nitrided surface regions formed at various processing conditions. They differ from the no-nitrided material mainly in the darkness of appearance, which can be referred to the presence of fine nitride precipitates. In the specimens processed at 470 oc, neither the compound layer nor the nitride network was found, Figs. 2,3. On the other hand, these structural constituents are clearly visible in the material processed at 500 o C for 60 min. and 530 o C for 120 min., respectively, Figs. 4 and 5. On the micrographs in Figs. 2-4, there is also the nitrogen diffusion depth marked out by yellow band on the right hand side. Note that there is not the nitrided region marked in Fig. 5 since it reaches deeper than the effective scale of the micrograph. Nitrides formed during the processing create individual formations embedded in the base material, Fig. 6. The size of the formations is of about several tens of nanometers. They have a lamellar structure with the thickness if individual lamellae ranging between 1 and 3 nm, Fig. 7. Previous investigations [10, 12] fixed that the nitrides are formed by iron and nitrogen and chromium and nitrogen, respectively nm 7 2 nm Figures 6, 7. HRTEM micrographs showing detail microstructure of nitrides in the region developed at 500 oc for 60 min. The saturation of the material with nitrogen increases the hardness, as shown in Fig. 8. Measurements revealed relatively high near-surface hardness even for the specimens treated at lower temperature and/or short dwell time. For the specimens processed at higher temperature and/or longer time, the near surface microhardness increases rapidly. The maximal values exceeded 1600 HV The main difference between specimens processed by various nitriding conditions is that the hardness after processing at lower temperature drops down at a shorter distance from the surface. On the other hand, only a slight hardness decrease was observed in the specimen processed at 530 o C for 120 min. These differences in microhardness depth profiles are also reflected in the surface hardness, which is markedly lower for the specimens processed at 470 o C (882 HV 10) than for those nitrided at 530 o C (1122 HV 10). Figure 9 demonstrates that the austenitizing temperature is a relevant factor influencing the three point bending strength, but only in the case of no nitrided material. In the case of the presence of the plasma nitrided layer it is clearly shown that this layer itself is the main factor influencing the toughness and the effect of the austenitizing temperature becomes much less significant. The toughness decreases again as the thickness of nitrided layer increases. The differences in the three point bending strength of no nitrided samples caused by different austenitizing temperature increase as the cross section of the specimens decreases,

113 In all of the specimens with no presence of a nitrided layer, the fracture surface exhibits clear symptoms of so - called low energetic transcrystralline fracture with a dimple morphology of the surface, Fig. 12a. As apparently shown, the initiation of fracture is realised through decohesion at the carbide matrix interfaces and fragmentation of carbides, also. These processes proceed mainly at the surface, in several centres and on the tensile strained side, as demonstrated in SEM micrograph in Fig. 12b. What is only one difference between the specimens austenitized at different temperatures (and having a different hardness) that from macroscopical point of view, the fracture surface of the specimen quenched from the lower temperature manifest a lot of secondary cracks, Fig. 12c. However, from microscopical point of view, these differences could only hardly be observed. The fact why the material austenitized at a lower temperature has higher fracture toughness can be considered as natural. It is known that the austenitic grain size increases as the temperature is raised and, also the products of the decomposition of the austenite (like martensite) can be expected to have coarser structure. These phenomena are well known as the limiting ones for the toughness and can clarify the lowering of the three point bending strength with increasing austenitizing temperature. a 100 m b 5 m c 750 m Figure 12. SEM micrograph showing the fracture surfaces of the no nitrided specimen, austenitized at: (a) 1050 o C - overview, (b) detail, (c) austenitized at 1000 o C The mechanism of fracture initiation in case of nitrided specimens differs clearly from that of no nitrided samples. The initiation is, in a way similar to the no nitrided material, located in a tensile strained face, but the fracture exhibits clearly the symptoms of transcrystalline cleavage, Figs. 13 a, b. As shown, the thickness of the cleavage region corresponds well with the nitrogen diffusion depth, see Table 1. Details of cleavage facettes from various places from the nitrided region are shown in Fig. 14. SEM micrograph in Fig. 14a shows the cleavage region at the surface, micrograph in Fig. 14b in the central area, respectively. Cleavage facettes manifest small steps that are connected

114 probably with the structure of the surface layer. Previous experiments clarified that the layer consists of the martensitic needles containing nitrogen, and ultra-fine nitride particles. Relatively coarse carbides are broken down during the propagation of the crack and they can also serve as the nuclei of the crack re initiation. a 5 m b 5 m Figure 13. SEM micrographs showing a cleavage fracture of the nitrided layer, developed at: (a) 470 o C for 30 min., (b) 530 o C for 120 min. a 5 m b 3 m Figure 14. SEM micrographs showing details from Fig. 13b (a) close the surface, (b) approx. 25 m below the surface The investigations of the fracture behaviour show that the cleavage propagation of cracks is connected with only a negligible plastic deformation. All of the energy input into the material is spent only for the formation of two new surfaces. This is a difference, compared to the no nitrided material, where a low plastic deformation of the material was found throughout the specimens and, as a consequence, the three point strength was remarkable higher. The lowering of material toughness with an increasing nitriding temperature and/or time can be explained through the fact that the portion cleavage region from the total cross section area increases as the thickness of nitrided layer increases. Although it is clear that the presence of nitrided layer at the surface lowers the toughness of the material, i.e. it acts in an undesirable manner, nitriding will still be used in an industry. Previous experiments confirmed that the nitriding increases the fatigue life time of tools made from P/M ledeburitic steels [13]. Other important aspect is an improvement of wear resistance, or, in many cases, also the improvement of adhesion of thin PVD films formed on pre-nitrided surfaces [14-18]. But, the conditions of the nitriding should be chosen carefully, in order to minimize the lowering of the toughness of the material.

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